Thermal-hydraulic analysis of a VVER-1000 core in MSLB ...

Thermal-hydraulic analysis of a VVER-1000 core in MSLB conditions

Svetlomir Mitkov1, Ivan Spasov

2*, and Nikola Kolev

1

1 Reactor Physics Laboratory, Institute for Nuclear Research and Nuclear Energy – Bulgarian

Academy of Sciences, Bulgaria 2 Department of Thermal Power Engineering and Nuclear Power Engineering, Technical University of

Sofia, Bulgaria

Abstract. The objective of this paper is to analyze the ability of a VVER-

1000 core and its control system to cope with a hypothetical main steam

line break (MSLB) accident in case of multiple equipment failures. The

study involves the use of advanced 3D core calculation models

benchmarked and validated for reactivity accidents in preceding studies. A

MSLB core boundary condition problem is solved on a coarse (nodal)

mesh with the coupled COBAYA/CTF neutronic/thermal hydraulic codes.

The core thermal-hydraulic boundary conditions are obtained from a

preceding full-plant MSLB simulation. The assessment of the core safety

parameters is supplemented by a fine-mesh (sub-channel) thermal-

hydraulic analysis of the hottest assemblies with the CTF code using

information from the 3D nodal COBAYA/CTF calculations. Thirteen

variants of a pessimistic MSLB scenario are considered, each of them

assuming a number of equipment failures aggravated by eight control rods

stuck out of the core after scram at different locations in the overcooled

sector. The results (within the limitations of the adopted modeling

assumptions) show that the core safety parameters do not exceed the safety

limits in the simulated aggravated reactivity accidents.

1. Introduction

A major concern in a main steam line break (MSLB) reactivity accident is the risk of core

overheating. In the computational analysis of such accidents the safety parameters of

particular interest are the fuel centerline temperature, the departure from nucleate boiling

ratio (DNBR) and the fuel rod cladding temperature. The fuel temperature should be kept

well below the UO2 melting temperature which can significantly vary depending on the fuel

burnup. As a VVER-1000 core can contain once, twice, three or four times burnt fuel

assemblies, it is of practical interest to analyze the consequences of such an asymmetric

reactivity accident at the end of core life and for various combinations of equipment failure.

This paper presents results from the analysis of thirteen variants of a hypothetical

MSLB scenario, each of them assuming a number of equipment failures plus eight control

rods stuck out of the core after scram at different radial locations in the overcooled sector.

* Corresponding author: [email protected]

PEPM'2021E3S Web of Conferences 327, 01013 (2021) https://doi.org/10.1051/e3sconf/202132701013

© The Authors, published by EDP Sciences. This is an open access article distributed under the terms of the CreativeCommons Attribution License 4.0 (http://creativecommons.org/licenses/by/4.0/).

The main objective is to demonstrate the use of state-of-the-art 3D core calculation

models which have been benchmarked and partly validated for VVER-1000 reactivity

accidents in preceding studies. A specific objective is to make a step in the analysis of the

ability of a VVER-1000 core and its control system to cope with such an accident without

exceeding the safety limits. Uncertainty analysis is beyond the scope of this work.

2. Methodology

2.1 Accident scenario

The considered accident scenario is based on aggravated variants of the pessimistic

scenario in the OECD/NEA VVER-1000 MSLB benchmark (V1000CT-2) [1]. The task is

to solve a MSLB core boundary condition problem using coupled 3D neutronic/thermal

hydraulic codes, given the core thermal-hydraulic (TH) boundary conditions as obtained

from a full plant simulation. The core boundary conditions (BCs) are taken to be as

specified in the V1000CT-2 benchmark [1, 2]. The plant transient is initiated at hot full

power by a large guillotine type break of steam line #4 outside the containment, upstream

of the steam intercept valve. The reference core is a real Kozloduy-6, Cycle 8 three-year

batch core at 270.4 EFPD [1]. The core contains once, twice and three times burnt UO2 fuel

of initial enrichment 4.23 w/o and 4.4 w/o. The steam generator feed-water valve in the

faulted loop fails to close on MSLB signal and remains open. The main coolant pump

(MCP) in the faulted loop fails to trip on MSLB signal and all MCP remain in operation.

The steam generator in the faulted loop continues uncontrolled cooling till the complete

evaporation of the secondary water. A cooler sector is formed at the core inlet, with

overcooling of up to 80°C. During the transient eight peripheral control rod clusters (CR)

are assumed to remain stuck out of the core after scram, all of them in the overcooled

sector. Thirteen cases with different radial combinations of the stuck rods are to be

analyzed to assess the values of the core safety parameters.

2.2 Codes and methods

2.2.1 Full-core nodal calculation

Full-core coarse-mesh (nodal) simulations for each CR configuration were carried out with

the coupled COBAYA4/CTF neutronic/thermal-hydraulic codes.

COBAYA4 [3, 4] is a 3D multi-scale core physics code using transport-corrected multi-

group diffusion approximation. It is developed by the Universidad Politecnica de Madrid

and benchmarked for VVER-1000 calculations [5, 6] in the frame of the EU NURISP [7]

and NURESAFE [8] projects. At the nodal level the analytical coarse-mesh finite-

difference (CMFD) method [9] is used. The code has radial mesh refinement capability.

COBRA-TF (CTF) [10, 11] is a recent version of the COBRA-TF thermal-hydraulic

(TH) code which uses a two-fluid, three-field modeling approach and has sub-channel

capabilities.

The COBAYA4/CTF coupling method [12] for VVER-1000 is based on the MED

Coupling libraries in the Salome platform [13]. The coupling and the coupled models have

been tested for VVER MSLB in preceding studies [5, 14]. The modeling assumptions in the

coupled COBAYA4/CTF VVER-1000 calculation models are briefly summarized below:

� Coarse-mesh COBAYA 3D neutron kinetics with:

- 30 axial nodes in the heated region;

- 2 nodes in each axial reflector;


2

- 6 triangles per hexagon.

� Use of a realistic VVER-1000 multi-parameter two-group cross-section library

[15] for reactivity accident analysis which has been generated with the APOLLO2

code [16] and validated in the frame of the NURESAFE project [8];

� Use of time-dependent assembly-by-assembly MSLB thermal-hydraulic core

boundary conditions (inlet temperatures, inlet mass flow rates, and outlet

pressures) obtained from a full plant simulation involving a quasi-3D reactor

pressure vessel TH model [17];

� Coarse-mesh CTF thermal-hydraulic model with one channel per assembly and 30

axial nodes in the heated part;

� Fuel model with 9 radial rings in the fuel, one for the gas gap and one for the

cladding;

� Temperature-dependent fuel and cladding thermal-physical properties [18];

� The spacer grids are not explicitly modeled and are taken into account by the

vertical pressure loss coefficients;

� The gas gap conductance coefficient is taken constant, equal to 3070 W/m2K, as

estimated at average core burnup of 26.6 MWd/kgHM [1];

� Chen’s model of nucleate boiling [19] and the W-3 critical heat flux (CHF)

correlation [20] with non-uniform power distribution.

2.2.2 Sub-channel assembly calculations

Two variants of the transient having the most risky values of the core safety parameters

were selected for sub-channel TH calculations of the hottest assemblies. Such simulations

were carried out for the hottest assemblies and assemblies next to them so that the analysis

includes once, twice or three times burnt fuel. The main modeling assumptions were as

follows:

� Assembly thermal-hydraulic problems were solved with inlet/output TH BCs from

the plant system simulation, and assembly powers and axial power profiles as

obtained from the full-core nodal simulation with COBAYA4/CTF;

� Radial pin-power distribution taken such as at hot full power for the hottest

assemblies, and with artificially imposed 5% radial tilt for the considered adjacent

assembly of higher burnup;

� Coolant-centered radial spatial mesh with 660 sub-channels per assembly;

� Axial mesh with 30 nodes in the heated region;

� Fuel model with 9 radial rings in the fuel pellet, 1 for the gap and 1 for the

cladding. The central hole is taken into account, and conduction in radial and axial

direction is considered;

� The bypass of 2.2% through the control rod guide tubes in the un-rodded

assemblies is not explicitly modeled and is taken into account by decreasing the

active coolant flow;

� The spacer grids are not explicitly modeled and are taken into account by the

vertical pressure loss coefficients;

� User defined coolant mixing coefficient of 0.01 (and exploratory option of 0.03 to

study the impact of higher coefficients);

� Use of the W-3 CHF correlation with non-uniform power distribution;

� Use of both Chen’s [19] and Thom’s [21] models of nucleate boiling, to compare

the performance;

� Constant fuel gap conductance coefficient equal to 3070 W/m2K as estimated at

average core burnup of 26.6 MWd/kgHM [1];

� Temperature-dependent fuel and cladding thermal-physical properties [18].


3

Details of the sub-channel input model for VVER assemblies can be found in ref. [22].

3. Results

3.1 Full-core nodal calculations

Table 1 Summary of the coarse-mesh computed core configurations and safety parameters

Case # Stuck CR, #

Max

return

to

power

MW

Hottest

assembly

# &

Adjacent

assembly

#

Assem-

bly

burnup

MWd/

kgHM

Fxy

Max

Tfuel,

ºC

Tmelt

UO2,

ºC

Max

Тclad,

ºC

Min

DNBR

1

90,91,105,

106,117,

118,130, 140

1098 104 / 117 15.3/ 31.0 7.694/ 6.958

2015/ 2018

2745/ 2650

335.5 4.21/ 4.27

2 82,90,91, 105,106,

117,118,130

889.8 104 / 117 15.3 8.927/

7.470

1851/

1580

313.2/

298.8

4.69/

5.98

3 79,90,91, 105,106,

117,118, 130

986.8 104 / 90 15.3 8.303/

6.964

1976/

1661

319 /

303

4.23/

5.57

4

63,90,91,

105,106, 117,118, 130

991.2 104 / 90 15.3 8.421/

7.440

1912 /

1680

319.2/

307.4

4.31/

5.29

5

63,82,90,

91,105, 106,117, 118

891.9 104 / 90 15.3/

24.1

8.986/

8.076

1826 /

1630

313.4/

303.7

4.72/

5.60

6

105,106,

117,118, 130,140,

142,151

1051 129 / 140 15.45 /

24.2 8.013/ 7.605

1944 / 1831

2745/ 2690

320.2 / 315.2

4.31/ 4.71

7

64,90,91,

105,106, 117,118, 130

933.5 104 / 117 15.3 8.703/

7.219

1902 /

1609

316.3 /

301

4.46/

5.80

8

90,91,94,

105,106, 117,118,130

957.3 104 / 117 15.3 8.388/

7.036

1912 /

1642

316 /

302

4.46/

5.69

9

90,91,105,

106,117,

118,120, 130

984.4 104 / 90 15.3 8.112/ 6.704

1926 / 1584

316./ 299.

4.43/ 5.96

10

91,105,

106,117,

118,120, 130,140

1005 129 / 140 15.45 8.141/

6.887

1957/

1665

318.8 /

304

4.27/

5.53

11

91,94,105,

106,117,

118,130, 140

963.7 129 / 117 15.45 8.301/ 6.989

1882 / 1636

315.4 / 301.5

4.54/ 5.70

12

91,105,

106,117,

118,130, 140,151

1035 129 / 140 15.45 8.115/

7.319

1906 /

1702

320 /

310.

4.30 /

5.11

13

91,105,

106,117, 118,130,

140,142

955.9 129 / 140 15.45 8.503/ 7.293

1884/ 1623

316.4 / 302.7

4.48/ 5.67

Limiting values 0.0 2840 1200 1.45*

of the safety parameters 49.0 2540 1200 1.45*

* When using the W-3 DNB correlation


4

Table 1 summarizes the suite of 13 variants of the MSLB scenario and the corresponding

coarse-mesh computed core safety parameters. In each case, the parameters of the hottest

assembly (which appears to be once burnt) and those of an adjacent assembly (twice or

three times burnt) with high Fxy power peaking factor are shown.

The nodal simulation results show that the highest return to power after scram and the

highest radial peaking factors occur in Case 1 and Case 6.

Figure 1 shows the time history of the total core power in Case 1 and Case 6 of the

transient. The corresponding maximum return to power after scram is 1098 MW,

respectively 1051 MW, at 67s from the onset of the transient.

Figure 2 illustrates the computed relative assembly powers at time of max return to

power (67s) in Case #1. The radial locations of the fully inserted control rod clusters are

marked in blue and those of the stuck rods – in beige. The hottest assembly is the unrodded

#104, marked in brick.

Figure 3 illustrates the computed relative assembly powers at time of max return to

power (67s) in Case #6. The core map format is as in Figure 2. The hottest assembly is the

unrodded #129, marked in brick.

The assembly-wise power distributions show that a big part of the core power is

released in the region of stuck rods in the overcooled core sector. The peak of the 3D power

distribution in the disturbed sector is near the core periphery. Correspondingly, Case 1,

assembly #104 (15.3 MWd/kgHM) and the adjacent #117 (31.0 MWd/kgHM), as well as

Case 6, assembly #129 (15.45 MWd/kgHM) and the adjacent #140 (24.1 MWd/kgHM)

were selected for additional sub-channel thermal-hydraulic analysis.

Figure 4 illustrates the coarse-mesh computed axial power distributions at time of max

return to power - core-averaged and assembly-wise for the assemblies of interest for sub-

channel analysis.

0

500

1000

1500

2000

2500

3000

0 30 60 90 120 150 180 210

To

tal p

ow

er,

M

W

Time, s

Case #1

Case #6

Fig. 1 Time history of the total core power


5

Fig. 2 Case #1 : Computed relative assembly powers at time of max return to power (67s)

Fig. 3 Case #6: Computed relative assembly powers at time of max return to power (67s)


6

0.5

0.6

0.7

0.8

0.9

1

1.1

1.2

1.3

1.4

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5

No

rmali

zed

axia

l p

ow

er

dis

trib

uti

on

Elevation, [m]

Case #1: Core averagedCase #1: Assembly#104Case #1: Assembly#117Case #6: Core averagedCase #6: Assembly#129Case #6: Assembly#140

Fig. 4 Computed axial power distribution at time of max return to power

3.2 Sub-channel thermal-hydraulic calculations

In Case 1, sub-channel TH analysis was carried out for the hottest assembly which is #104

(15.3 MWd/kgHM) and for the adjacent #117 (31.0 MWd/kgHM).

In Case 6, sub-channel calculation was performed for the hottest #129 assembly (15.45

MWd/kgHM) and the assembly #140 (24.1 MWd/kgHM) next to it.

In both cases the assembly inlet and outlet TH boundary conditions at time of max

return to power are nearly the same. The inlet temperatures for Case 1, assembly #104 and

#117 are 207.62 ºC, and for Case 6, assembly #129 and #140 are 207.7 ºC and 207.66 ºC

respectively. The corresponding inlet mass flow rates are nearly the same.

It should be noted that some preliminary sub-channel results for Case 1, hot assembly

#104 have been reported in ref. [23], and this study extends the sub-channel analysis to

more cases and various assemblies including such of higher burnup, where the UO2

melting temperature is lower.

The peak of the 3D power distribution in the disturbed sector is near the core periphery

and corresponds to the radial location of the hottest assembly. For such a configuration and

in the lack of full-core pin-by-pin calculation, it is reasonable to assume a hot full power-

like radial pin-power distribution for the hottest assembly (#104 in Case 1, and #129 in

Case 6). An approximate tilt of ±5% across the assembly section is assumed for the

adjacent assemblies.

Figure 5 illustrates the adopted pin-power distribution at hot full power (HFP).

Details of the computed distributions are illustrated in Figures 6 through 14.

Table 2 summarizes the sub-channel results for the core safety parameters, obtained

with the Chen and Thom options of the nucleate boiling model in CTF. The results in this

table and the figures below (except for Figure 12) are obtained with user defined cross-

channel mixing coefficient of 0.01. A small value is specified because this coefficient is not

yet validated. Thus, the cross-channel mixing is mainly gradient driven. See Figures 11 and

12 for a parametric study of the impact of higher values of the user defined mixing

coefficient on the vapor volume fraction.


7

0.000

1.001 1.001

1.000 1.000 1.000

0.998 0.997 0.000 0.998

0.992 0.984 0.993 0.994 0.992

0.000 0.990 0.992 0.993 0.992 0.000

0.982 0.986 0.988 0.000 0.988 0.987 0.983

0.971 0.973 0.975 0.985 0.985 0.975 0.973 0.971

0.985 0.978 0.976 0.979 0.979 0.979 0.976 0.978 0.985

1.015 1.002 0.998 0.993 0.995 0.995 0.994 0.998 1.002 1.015

1.078 1.057 1.046 1.041 1.040 1.038 1.040 1.054 1.046 1.057 1.078

Fig. 5 PERMAK [24] computed radial pin power distribution in a fuel assembly at hot full power

Table 2 Sub-channel computed core safety parameters

Case Assembly #

/ Burnup,

MWd/kgHM

Fxy 5% tilt of

the radial

pin-power

distribution

Hot cell

exit vapor

volume

fraction,%

Max

T

fuel,

ºC

T

melt

UO2,

ºC

Max

Тclad,

ºC

Min

DNBR

1 104 / 15.3 7.733 no 0.32 2506 2745 333.1 2.66

1 117 / 31.0 6.996 yes 0.20 2367 2650 328.3 2.96

6 129 / 15.45 8.013 no 0.30 2456 2744 332.7 2.69

6 140 / 24.2 7.605 yes 0.31 2439 2690 332.4 2.71

6 140 / 24.2 7.605 no 0.16 2341 2690 327 3.03

6* 129 / 15.45 8.013 no 0.12 2456 2744 330.1 2.69

6* 140 / 24.2 7.605 no 0.05 2341 2690 327 3.03

*When using the Thom nucleate boiling correlation

In Table 2, the limiting values of the core safety parameters are as follows:

� Tmelt UO2, ºC: 2840 ºC for fresh fuel and 2540 ºC for burnt fuel (at 50 EFPD for the

considered VVER fuel);

� Tclad: 1200 ºC;

� Min DNBR: 1.45 (when using the W-3 DNB correlation).

The predicted values of the core safety parameters (within the limitations of the

modeling assumptions) do not exceed the safety limits. However, the impact of the 5% tilt

of the radial assembly pin power distribution is considerable and indicates that this issue

requires further attention using full-core pin-by-pin coupled calculations.

Figures 6 and 7 show the sub-channel results for Case 1, assembly #104 obtained with

HFP pin-power distribution and Chen’s nucleate boiling model.

Figure 8 illustrates the results for Case 1, assembly #117, obtained with 5% tilt in the

pin-power distribution across the hexagonal section and Chen’s nucleate boiling model.

Figures 9 through 12 show the results for Case 6, hot assembly #129 (15.45

MWd/kgHM). The results аrе obtained using HFP pin-power distribution. The comparison

of the sub-channel liquid fraction maps in Figures 11 and 12 shows the impact of using

higher values of the cross-channel mixing coefficient (0.03) – the sub-cooled boiling area

shrinks toward the assembly center.

Figures 13 and 14 show the results for Case 6, assembly #140 (24.2 MWd/kgHM),

obtained with ±5% tilt in the pin-power distribution across the hexagon and Chen’s

nucleate boiling model.


8

The results in Figures 6, 8, 9 and 13 show that despite the high power peaking factors in

the hottest assemblies only subcooled boiling and small vapor volume fraction in the upper

part of the assemblies is predicted, because of the low inlet temperatures.

Figure 9 shows that the use of Chen’s nucleate boiling model predicts higher vapor void

fraction in the subcooled region compared to that obtained with the Thom correlation.

Both Chen’s and Thom’s correlations predict nearly the same axial profiles of the fuel

centerline and cladding temperatures (see Figure 10).

1200

1400

1600

1800

2000

2200

2400

2600

2800

0

0.0005

0.001

0.0015

0.002

0.0025

0.003

0.0035

0.004

0 0.5 1 1.5 2 2.5 3 3.5

Tem

pera

ture

, [C

]

Vap

ou

r fr

acti

on

, [-

]

Elevation, [m]

Fig. 6 Case #1, Assembly #104: Computed hottest pin centerline temperature and hot cell

vapor volume fraction at max return to power, when using the Chen nucleate boiling model

0

1

2

3

4

5

6

7

8

9

0 0.5 1 1.5 2 2.5 3 3.5

DN

BR

, [-

]

Elevation, [m]

Fig. 7 Case #1, Assembly #104: Computed DNBR (z) using the W-3 correlation and HFP pin

power distribution, at max return to power


9

1000

1200

1400

1600

1800

2000

2200

2400

2600

2800

3000

0

0.0002

0.0004

0.0006

0.0008

0.001

0.0012

0.0014

0.0016

0.0018

0.002

0 0.5 1 1.5 2 2.5 3 3.5

Tem

pera

ture

, [C

]

Vap

ou

r fr

acti

on

, [-

]

Elevation, [m]

Fig. 8 Case #1, Assembly #117: Computed hottest pin centerline temperature and hot cell

vapor volume fraction at max return to power, when using the Chen nucleate boiling model

0

0.0005

0.001

0.0015

0.002

0.0025

0.003

0.0035

0 0.5 1 1.5 2 2.5 3 3.5

Vap

ou

r fr

acti

on

, [-

]

Elevation, [m]

Thom's correlation: Bundle averaged

Thom's correlation: Hottest sub-channel

Chen's correlation: Bundle averaged

Chen's correlation: Hottest sub-channel

Fig. 9 Case #6, Assembly #129: Computed axial distribution of the bundle averaged and hot cell

vapour volume fractions


10

200

220

240

260

280

300

320

340

360

1200

1400

1600

1800

2000

2200

2400

2600

2800

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5

Tem

pera

ture

, [C

]

Tem

pera

ture

, [C

]

Elevation, [m]

Thom's correlation: Tfuel

Chen's correlation: Tfuel

Thom's correlation: Tcladding

Chen's correlation: Tcladding

Fig. 10 Case #6, Assembly #129 : Computed hottest pin centerline and cladding temperatures.

Tmax, fuel=2456,4 C , Tmax,clad-Chen = 332,7 ºC, Tmax,clad-Thom = 330,6 ºC

Fig. 11 Case #6, Assembly #129 at max return to power: Computed exit liquid volume fraction, with

user defined mixing coefficient of 0.01 and Thom’s’nucleate boiling model


11

Fig. 12 Case #6, Assembly #129 at max return to power: Computed exit liquid volume fraction, with

user defined mixing coefficient of 0.03, and Thom’s nucleate boiling model

0

0.0002

0.0004

0.0006

0.0008

0.001

0.0012

0.0014

0.0016

0.0018

0.00 0.50 1.00 1.50 2.00 2.50 3.00 3.50

Vap

ou

r fr

acti

on

, [-

]

Elevation, [m]

Thom's correlation - hottest cell

Thom's correlation - bundle averaged

Chen's correlation - hottest cell

Chen's correlation - bundle averaged

Fig. 13 Case #6, Assembly #140: Computed vapor volume fraction with pin-power distribution tilt


12

220

240

260

280

300

320

340

0

500

1000

1500

2000

2500

3000

0 0.5 1 1.5 2 2.5 3 3.5

Tem

pera

ture

, [C

]

Tem

pera

ture

, [C

]

Elevation, [m]

Tfuel

Tcladding

Fig. 14 Case #6, Assembly #140 : Computed hottest pin centerline and cladding temperatures using

Chen’s nucleate boiling model and pin-power tilt. Tmax,fuel = 2439 ºC, Tmax,clad = 332.4 ºC

4. Summary and conclusions

The safety parameters of a VVER-1000 core near the end of life were analyzed for a suite

of pessimistic MSLB accident scenarios. For this purpose, coupled full-core

COBAYA4/CTF neutronic/ thermal-hydraulic simulations at the nodal level were

supplemented by CTF sub-channel calculations for the hottest assemblies of different

burnup in the most risky cases.

In the CTF sub-channel assembly simulations, modeling options in the modeling

assumptions with the Chen and Thom nucleate boiling models and with different user-

specifies cross-channel mixing coefficients were compared.

The results show that:

� In all cases the predicted fuel cladding temperature at time of max return to power

is far below the limiting value of 1200 ºC;

� The predicted sub-cooled boiling in the upper part of the hottest sub-channels is

insignificant, even in the worst cases. The Chen model of nucleate boiling predicts

higher vapor volume fraction compared to that obtained with Thom’s model (as

observed in other LWR assembly calculations);

� The minimum DNBR value of 2.66 predicted in the worst Case 1, Assembly #104

is well above the limiting value of 1.45 when using the W-3 CHF correlation;

� The predicted fuel centerline temperature in the hottest assemblies has a moderate

margin of about 237 – 251 ºC to the burnup dependent UO2 melting temperature.

Because of its importance, and of its strong dependence on the local pin power

distribution and on the gap conductance coefficient, the assessment of this safety

parameter deserves further attention.

� In case of boiling in the assembly, the user-defined cross-channel mixing

coefficient in sub-channel CTF needs further validation on measured assembly


13

data because of its impact on the void fraction distribution. In the meantime, small

values or not using this coefficient produce more conservative results.

The results (within the limitations of the adopted modeling assumptions) show that the

core safety parameters do not exceed the safety limits in the simulated aggravated reactivity

accidents.

Further improved analyses are expected to include a coupled pin-by-pin

neutronic/thermal-hydraulic calculation, at least for a mini-core of seven assemblies around

the hottest one, as well as a dynamic gap conductance coefficient, a DNBR from VVER-

specific CHF skeleton tables in CTF, and uncertainty quantification.

5. References

1. N.P. Kolev et al: VVER-1000 Coolant transient benchmark Phase II (V1000CT-2),

Vol.2, Final Specifications of the VVER-1000 MSLB problem. NEA/NSC/DOC

2006(6) © OECD (2010)

2. N.P. Kolev, I. Spasov, Tz. Tzanov, E. Royer: VVER-1000 Coolant Transient

Benchmark: Phase 2 (V1000CT-2) Volume 4: Summary results of coupled 3D

kinetics/core-vessel thermal hydraulics and core-plant MSLB simulation.

NEA/NSC/DOC (2011)3, Paris © OECD (2011)

3. COBAYA team: COBAYA4 user’s guide. UPM Report, Madrid, (2015)

4. C. Ahnert: Capacities and achievements of the COBAYA4 code after the NURESAFE

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